Composite connections in steel and concrete. Part 2 — Moment capacity of end plate beam to column connections

0143-974X(95)00015-1

J. Construct. Steel Res. Vol. 37, No. 1, pp. 63-90, 1996 Copyright © 1996 Elsevier Science Ltd Printed in Great Britain. All rights reserved 0143-974X/96 $15.00 + 0.00

ELSEVIER

Composite Connections in Steel and Concrete. Part 2 Moment Capacity of End Plate Beam to Column Connections •

a ~

Y. X l a o , ' B. S. C h o o b & D. A. Nethercotb "Department of Civil & Environmental Engineering, University of Southampton, Southampton, SO17 1BJ, UK bDepartmem of Civil Engineering, University of Nottingham, Nottingham NG7 2RD, UK (Received 11 July 1994; revised version received 29 March 1995; accepted 16 May 1995)

ABSTRACT A comprehensive mathematical model has been derived to predict the behaviour of different types of composite connection. Its derivation for the prediction of moment capacity for composite endplate type connections is presented herein. The model has been validated using available test results from the database developed by the authors. A parametric study has been carried out using this model to check the balance between connection and beam strength for a wide range C~ variables. Suggestions are made for composite endplate connections to be rationally designed as partial strength, moment resisting connections.

INTRODUCTION It is now more than 20 years since the first studies' of the effects of composite action on the performance of beam to column connections were undertaken. Since then numerous sets of experiments 2 have been conducted in various centres in an attempt to understand the role of the different variables on the performance of composite connections. More recently, systematic studies of the types of connections popular in multi-storey steel frames, designed on a n o n - s w a y basis 3"4 and incorporating metal deck composite flooring, have been undertaken. 5 Results from these tests have been drawn together into a substantial compaterised database 6 that has permitted the systematic study of the effects of changing individual parameters. As a result, it has now become

*Previous address: University of Nottingham, UK.

64

Y. Xiao, B. S. Choo, D. A. Nethercot

possible to derive design methods based on the principles of the force transfer system operating within the connections that is provided by the combination of the steelwork detail, the shear studs and the slab reinforcement. It is the purpose of this paper to present such a design method for the particular case of composite endplate beam to column connections. The approach adopted follows the basic design philosophy of both BS 5950: Part 3.17 and EC4, s although neither of these documents actually present design rules for composite connections. Where necessary, reference to the associated detailed rules for the steelwork part, e.g. bolt strengths, is made to suitable national and international steelwork codes. 3,4 Thus, the approach used is based on the concept of stress blocks as the model for determining the contribution to the overall moment capacity of the connection from the individual elements. The specific procedures for predicting the moment capacity of composite endplate will be described herein. For cleated and tinplate arrangements, 5 although an essentially similar model has been adopted, different detailed procedures are required. The derivation and validation of these procedures will be reported elsewhere. The importance of being able to accurately predict the moment capacity of composite connections as one of the key elements in a design method for non-sway composite frames has been discussed elsewhere.9 In the quasi-plastic frame design approach proposed, four key measures of behaviour associated with connection performance were identified as: connection moment capacity, connection stiffness, connection rotation capacity and rotation requirement in the connections. The last of these has been fully investigated elsewhere, both at a fundamental level ~° and in terms of a more simplified approach that links the necessary predictions directly to the available degree of moment redistribution in the system. 11 Prediction of connection rotation capacity has also been considered in Ref. 12 using an approach that is fully compatible with that presented herein to cover strength. DESIGN APPROACH Basis

The derivation of formulae to predict the moment capacity of composite endplate beam to column connections presented herein is based on the following assumptions: • Any tensile strength remaining in the concrete slab after cracking is neglected. • All reinforcement in the slab reaches yield, but any strain hardening is neglected.

65

Composite connections in steel and concrete

• In the compression zone of the connection, yield will be attained and the resulting distribution of forces will be uniform. Each of these is based on detailed observations of behaviour in the authors' own tests, 5 as well as an examination of the background papers and reports used to produce the database. 6 The concrete slab after cracking is incapable of transferring any tensile force except through the rebars and other possible forms of reinforcement (such as metal decking). Through measurement of the strains in the reinforcement it is known that extensive yielding of the reinforceme,nt normally occurred across the whole slab section. Strain measurements also detected that plasticity was present in the region subjected to compression. Thus, the above follow directly from the behaviour observed in tests and simplify the derivation. The ability of the metal decking to contribute tensile resistance is included only when the decking is oriented parallel to the direction of the beam. This follows test evidence that metal decking arranged in this way can be treated as partial replacement for rebar reinforcement. 5 Neglecting (for the present) any tensile contribution from the bolts in the endplate, the basic system of force transfers assumed to develop the connection's moment capacity is shown in Fig. 1. This was the model originally proposed in Ref. 13, which did not include an), contribution from the bolts and which assumed the lever arm to be fixed as the distance from the reinforcement to the beam bottom flange. From the tests, it is known that the neutral axis is usually positioned below the first row of bolts. Thus, the simplification of Fig. 1 is too conservative and cannot accommodate changes to the steel details. Ideally, any bolt in tension (above the neutral axis) should be considered in the establishment of the moment model. Resistance to vertical shear, i.e. transfer of the beam's end reaction into the column, is assumed to be provided solely by the steel part of the connection. Although the slab may make some contribution due to friction between the

/'--X T/--X

--FY;A } Mp

Mp = A Sfy df

% Fig. 1. Simple load transfer in a composite connection.

66

Y. Xiao, B. S. Choo, D. A. Nethercot

cracked surfaces and dowel action of the longitudinal reinforcement, this has not yet been studied fully. Test data on the interaction of shear and moment in composite connections are sparse and somewhat inconclusive; there does not, however, appear to be any significant reduction in moment capacity due to the presence of a high coincident shear. Thus, shear capacity should be determined using procedures for the equivalent bare steel joint. H The basic steps in determining the connection's moment capacity using the concept of a rigid-plastic, stress block approach consistent with the accepted methods for calculating the moment capacity of composite beams 7,8 are therefore with reference to Figs 2 and 3: (i) (ii) (iii) (iv)

Determine tensile force Fs based on yield of the reinforcement. Calculate the available compressive capacity Fb. Determine the position of the neutral axis. Decide whether any additional tensile resistance Ft is required from the bolts. (v) Make any adjustments to the position of the neutral axis and the force components. (vi) Locate the centroid of either the total tensile force or compressive force, determine the lever arm and hence calculate the moment resistance Mu.

Tensile resistance from composite floor Tensile force, Fy in Fig. 1, is provided by the longitudinal slab reinforcement (plus the decking if its ribs run parallel to the steel beam). It is assumed that

Fig, 2. Internal force diagram of beam to column model.

Composite connections in steel and concrete

67

Calculate tensile force

Fs

Determinecompressiveforce

Fb

I

Locateneutralaxisposition

,

Locatecen~roid ~ pZosition[ _

~-~

b

I Determine,F/TtIt tensilef°reeI

Ultimatemomentcapacity Mu

Fig. 3. Frame chart for procedure of moment capacity prediction.

sufficient shear studs are present for full interaction7,8 to occur, i.e. F~ is not controlled by shear failure at the steel-concrete interface. However, design on the basis of partial interaction is acceptable in the hogging moment region even though there is a reduction of the moment capacity for the connection. In certain circumstances the benefit of slip of the shear studs may be helpful as a contribution towards increasing the rotation capacity. This point is addressed elsewhere.J5 The following steps give the procedure for calculating the tensile force provided by the concrete floor. The design tensile force of the reinforcement in the slab is:

or

Fy=:O.87fyA~ Fy=O.87LA~+O.91LmAm

if metal de,cking is parallel to the beam direction where f,, = Characteristic strength of rebar A~ = Section area of reinforcing bars fym = Characteristic strength of metal decking Am = Section area of metal decking.

(1) (2)

68

Y. Xiao, B. S. Choo, D. A. Nethercot

The design resistance of shear connectors (example of headed shear studs) within the hogging moment region is: Q = n Q~

(3)

where n = Shear studs number Q, = Design strength of shear studs. The calculation for the shear strength of the shear studs will be different depending on the specific code used. If the concrete slab is cast on metal decking the resistance of the shear connectors should be reduced by multiplying by an appropriate factor k which can be obtained from the relevant code. 7'8 The tensile force F~ provided by the concrete slab will be depended upon the following conditions: If F y > nQ,,.

The tensile force of the concrete floor is: Fs = n Q,,.

(4)

If Fy <- n Qk .

The tensile force of the concrete floor is: = G-

(5)

Compressive resistance The minimum compressive force that could be transmitted corresponds simply to the compressive resistance of the whole of the area of the steel beam (or a proportion thereof if the beam cross-section classification 3'4 is such that it is only partly effective). It may therefore be obtained from: (6)

Fb =pyAb

where py Ab If the beam section should

= Design strength of steel member = Beam section area. section is semi-compact or its web is slender, the effective be used for calculating strength (Fig. 4).

Fb =py(Ab - (d - d')tbw)

(7)

Composite connections in steel and concrete be

~ I

69

compression [ tension

o.s7 f,

I,

[[

I

P.N.A -- -- --~..~..tNeglcct

Hx

,

I

V--

1119 te

Py Iz= (275/I:h,)vz

Note. P.N.A denotes plastic neutral axis of effective section Fig. 4. The effective section of class 3 semi-compact section from Ref. 7.

where d = Depth of the beam web between the fillets d' = Depth of beam web excluded from calculation tbw = Thickness of beam web. If the beam section has semi-compact flanges or a slender web, only the elastic strength may be used in formula 6, which should then be referred to Ref. 7.

Moment capacity Knowledge of the potential tensile and compressive resistances available from the slab and the beam permits the location of the neutral axis to proceed. Four cases are possible: (i) (ii) (iii) (iv)

Neutral Neutral Neutral Neutral

axis axis axis axis

is is is is

within within within within

the the the the

concrete slab beam's upper flange beam's web above the top row of bolts beam's web below the top row of bolts.

For the normal range of reinforcement ratios (e.g. from 0.7% to 1.4%), the first three cases are much less likely compared with case (iv), because these cases are usually associated with very high ratios of reinforcement for flush endplate connections. The first case, however, is very likely for partial depth endplate connections, especially for when the endplate is welded to the beam web only. For case (iv) some rows of bolts will be in tension and will therefore augment tlae tensile resistance determined from eqns (4) or (5).

Y. Xiao, B. S. Choo, D. A. Nethercot

70

Case (i) This case is associated with the condition shown in Fig. 5(a) and defined by

Since the available compressive resistance is less than the available tensile resistance, part of the slab is required to transfer compression. Clearly this must be given by: Fc = Fs -

(8)

and the depth of concrete in compression Xp may be obtained from:

Xp = 0.45(0.55befJ

(9)

where be = Effective width of composite slab fen = Concrete characteristic cube strength. The concrete is assumed to be stressed to a uniform compression of 0.45fcu over the full depth of concrete in compression. A factor 0.55 has been included to allow for the case where the ribs of the decking are parallel to the steam beam; for decking spanning transversely this factor should be taken as unity (Fig. 6). If a full depth of concrete does not intersect with the column face then the calculation will shift to case two as no concrete will be in compression. It is recommended that, where possible, trough with a full concrete section should intersect with the column face to improve the bond between the concrete and the rebar. Knowing the position of the neutral axis, the connection moment capacity Mo may be determined by taking moments about the centroid of the reinforcement:

(lO)

where D dc

and d c

= Depth of steel beam = d - c - ~b/2 = Depth of slab = Thickness of concrete cover

Composite connections in steel and concrete

71

b~

_t I~ d

--I

~

m

axis ====================:==~ : : : , ~

°I__

II',1 "t" II "r

.

J_ V ~

..........

=========================

-

tbf

~ I

f~

I_

-..

F$ d~- x p f xp .

~r~-

.

Fb

Fb

Case ii (b)

Casei Ca)

F$

Fs

Bolt ~ w 1 Bolt ~ w i Bolt row k

--~

--~1T
p 2/3p.

Fb "~:

dI

Fb

~.~

Case iii

Case iv

(C)

(d)

Fig. 5. Different positions of neutral axis in the concrete section.

Y. Xiao, B. S. Choo, D. A. Nethercot

72

Xp' Xp'- 0.55Xp

(a)

(b)

Ribs run transverse to beam

Ribs run parallel to beam

Fig. 6. Effective depth of concrete section in compression.

~b

= Diameter of the rebar.

C a s e (ii)

This case is associated with the conditions shown in Fig. 5(b) and defined by: F,
but Fb -- Fs -< P y f l tbf where f~ = Width of beam flange tbf = Thickness of beam flange. The depth of beam flange not in compression Xp may be obtained from: Fs - Fb xp-

(1 1)

Pyfl

The distance of the centroid of compression from the bottom surface of the steel beam dzl is:

(tbf -- Xp)f~(D a,, =

1 - ~tbe --

1

~Xp)

1 + 2-(D -

2tbf)tbwD +

(tbf -- Xp)f~ + ( D - 2tbe)tbw + f~tbe

1 ~f~tb 2

(12)

and the moment capacity Mu can then be obtained by taking moments about the centroid of the compressive area to give: Mu = Fs(dc + D - dzO .

(13)

Composite connections in steel and concrete

73

Case (iii)

This case is, associated with the condition shown in Fig. 5(c) and defined by: Fb -

F s :> Pyfltbf

but Fb -- F~ Pyfltbf ~ Py(Pl -- tbf)tbwin which p~ = distance from top bolt row to top surface of the steel beam. The depth of beam web not in compression Xp is given by: -

--

Xp =

-

PyfAf

(14)

tbwPy

and the position of the centroid of compression is defined by: tbw(D - 2 t b f - - Xp)(D - Xp) + f l t b f 2 dz2 =

2(tbw(D -- 2tbf -- Xp) +fltbf)

(15)

The moment capacity Mu may then be obtained from: Mu = C(d~ + D - dz2 ) .

(16)

Case (iv)

This case is associated withthe condition shown in Fig. 5(d) and is defined by: Xp>Pl.

Some bolts rows will now be in tension and the neutral axis must be located by checking to see how many rows need to be in tension to produce longitudinal equilibrium. In this case the bolt rows in the tension zone will all be assumed to yield except for the lowest row. This assumption has been carefully examined by looking in detail at results from several tests and analyses. Tests using special bolts (with strain gauges) have been conducted to check the tensile strains in the bolts/6 The results indicated that all bolt rows in the tension zone will reach yield. Numerical analysis for selected test specimens using ABAQUS has also shown that both the bolts and endplate in the tension region reach yield at the ultimate stage./7 From the combination of tests, numerical and theoretical work, it is therefore considered that the assumption is fully justified. With the presence of a suitable amount of reinforcement in the concrete

Y. Xiao, B. S. Choo, D. A. Nethercot

74

slab the tension bolts will not be overstressed at the ultimate stage, although they may reach yield (bolts may reach yield first). Plasticity will develop in the tension region. For all the composite endplate connection tests (more than 50) gathered by the authors fracture of the bolts has only occurred in those of Ref. 18. Three tests were reported in which the top row of bolts fractured. These cases were associated with comparatively thick endplates (15 mm), small bolts (19 mm diameter) and low level of reinforcement (0.74%). To improve the design of composite connections and to eliminate brittle failure modes specific measures have been proposed in Refs 5, 15. Bolt tensile strength will actually be governed by the bolts and their connecting components such as: endplate, column flange and welds, etc. dependent on which yields first. This has been fully discussed in Refs 15, 19. Yielding of the bolts is not the sole governing factor in the tension zone. The appropriate calculation models can be found in Refs 3, 15, 20. Individual bolt strength Ft has been used to summarise the governing conditions of both the bolts and their connecting components. Assume k rows of bolts to be in tension. A region extending to two thirds of the vertical pitch directly below the lowest row of bolts is assumed to be in tension. This assumption will be validated in the later part of the paper. Internal force equilibrium yields the condition: 1 F~ + 2Ft k = pyfltbf + p y t b w ( D -- Pl -- tbf -- (k -- ~)p)

(17)

where p Ft

= Bolt pitch = Individual bolt tensile strength (governed by bolts and their connecting parts) which may be arranged to give: 1 p y f t b f - Fs + p y t b w ( D -- Pl -- t b f + ~ p )

k=

2Ft

+

pytbwP

(18)

If k is smaller than 1, one row of bolts will be treated as being in the tension area and the tensile strength of a single bolt should be multiplied by k. A bolt row will reach yield when k is 1. The neutral axis could then be conservatively relocated to the centre line of the second row bolts (as discussed in the later part of the paper). If k is greater than 1, the tensile bolt row number k, which has been obtained from the formula, will be a real number where the integer part i represents the rows of bolts which have yielded, and the decimal part

C o m p o s i t e c o n n e c t i o n s in steel a n d concrete

75

j represent,; the stress factor for the row of bolts which has not yielded and is adjacent to the neutral axis. For either case writing k as: (19)

k=i+j

in which i = Number of rows of bolts reaching yield j -- Stress factor for the bolt row adjacent to the neutral axis. The total rows of bolts in the tension zone will be i + 1. For example, when k <-- 1 then i = 0 and k = j and total number of bolts in the tension zone is one row. Define the distance from the bottom surface of the steel beam to the centroid of compression as: tb~[(O

-- Pl -- ip --

2p) 2 -- tb2] +fltb 2 , , J

d~3 =

2 2[(0

-

p~ -

ip -

(20)

~p

-

tbOtbw

+fltbf]

Thus, the distance between the centre of the reinforcement and the centroid of the compressive area becomes: dc + D -- d~3

(21)

and the contribution to the connection's moment capacity from the reinforcement M~ is: Ms = F~(dc + O

-

C3)-

(22)

All the bolt rows in the tension region will also contribute to the moment capacity. Taking each row in turn gives: 2 M, = .,.Ft(D -

p, -

ip -

~p -

dz3)

M2 = 2Ft(D

-

p, -

(i -

I)p -

2 ~p -

Mi = 2F,(D

-

p, -

p -

2 jp

2 M i + , = ?-j F , ( D

-

p, -

~p -

-

dz3)

dz3)

dz3)

or summarising as a mathematical progression:

(23)

76

Y. Xiao, B. S. Choo, D. A. Nethercot : M, + M2 + . . .Mi + Mi+,

= 2Ft[k(D

- Pl -

i(i +

2 ~P -

dz3)

1)]~,

2

(24)

"

The moment capacity Mu for a composite connection can be derived by adding formulae (22) and (24) together Mu = Fs(dc + D +2F,[k(O

(25)

dz3)

- pl -

2 ~p -

dz3) -

i(i + 2

1)].p "

For the most common situation when only one row of bolts is in tension this reduces to: 2 M o = F s ( d ~ + D - dz3 ) "]- 2 j F t ( D

- P l - - ~ P - - dz3) .

(26)

Moment capacity with an unstiffened column web If the column web is unstiffened, the resistance of the column web in the compression region needs to be checked. The column web and flange are both potentially critical components. Three failure modes are possible: • Buckling of the web over most of the depth of the member. • Crushing of the web close to the flange, accompanied by plastic deformation of the flange. • Crippling of the web. These checks are covered by current steel codes 3"4 and there is thus no need to further describe the calculation procedures. However, the effective compression area of the column web needs to be defined for the buckling of the column web. As shown in Fig. 7 the loaded length from the endplate to the column web is

bolt 2 D

oo.

I/3p

***

bolt i+ 1

Fig. 7. Compressive length of column web under the Ilush endplate.

Composite connections in steel and concrete

2 s = [D - Pl - (i - ~p)].

77

(27)

The effective loaded width of the column web could be referred to Ref. 3 as: (28)

h e tf : (s 2 + h2) 1/2

where h = Depth of steel column. From the formulae (27) and (28) the effective buckling section area for the column web can be determined. The buckling strength Nb may then be obtained using Refs 3, 4. Usually this strength is smaller than the compressive strength of the beam end. Therefore, the compressive strength Fb will be: Fb = Nb,

(29)

Similarly, the crippling strength Nc and yielding strength Ny of the column flange can be obtained and substituted into eqn (29) as shown in Ref. 20. Comparison of eqn (27) with eqns (4) or (5) thus permits the governing condition to be identified and the governing strength of the connection to be calculated. Column buckling failure should normally be excluded from the design. The equivalent check for a partial depth endplate type of composite connection (partied depth endplate may be positioned at any position on the beam) is conducted in a similar way. This will not be discussed herein. The only significant difference is in the calculation of the compressive resistance of the beam, for 'which A b in Formula (6) is the part of the beam section area which is covered by the partial depth endplate.

VALIDATION

Full depth endplates Availability of the computerised database of composite connection test results 6 has permitted the design procedure outlined above to be validated by comparing predicted failure loads with the test values. Unfortunately not all results in the database are suitable for validation purposes due to the absence of key items of information in the test reports. This is a frequently encountered source of frustration when seeking to use reported test data. In certain other cases some of the more detailed aspects of the test arrangement w e r e not reported,

78

Y. Xiao, B. S. Choo, D. A. Nethercot

but it has been considered reasonable to use typical values, e.g. where concrete cover was not reported it has been taken as 25 mm, as a way of increasing the percentage of directly usable tests. Material data was not available for the tests of Ref. 18 and nominal values have been used for the calculations. Although predictions are on the safe side, the results have not been included in the final statistical summary. All the partial safety factors have been removed from the formulae to give direct comparison with test results. Forty-five tests have been used for validation including the flush endplate and partial depth endplate connections. All four cases in the model have been cross-examined by the test results. The comparison of the calculation and test results is listed in Tables 1, 4. The possible location of the neutral axis is dependent on the joint detail and reinforcement ratio, etc. For all the tests of flush endplate connections examined, only for case three and four were test results available, with the majority being for case four which means some bolt rows will be in tension. Analysis using the model indicates that a very high percentage of reinforcement in the slab is required to produce case (i) and (ii) for composite flush endplate connections. It is, in fact, not very economical to design the composite connection in this way since the situation is similar to reinforced concrete members which have been over-reinforced. For the partial depth endplate, case (i) and (ii) test results are available, particularly for the case when the endplate was welded to the beam web only since the steel part could not then provide a high compressive strength. The neutral axis moves into the upper position of the connections, sometimes even into concrete slab which corresponds to case (i). Six different groups of tests from the UK, Italy and France, 5'~6.lS'2j-z3 comprising the 30 separate specimens listed in Table 1 have been used. All specimens were of a cruciform arrangement representing an internal joint and most used symmetrical, i.e. balanced loading. The connections have 2, 3, 4 bolt rows, respectively. Only the tests of Benussi 22 and Law 23 employed solid slabs. Table 2 lists the comparison between test results and predictions. Tests in which mesh fracture occurred are not included in the statistical summary. Tests from Ref. 18 have also not been included in the final statistical summary. The comparison shows most of the predictions to be on the safe side; when the prediction was higher than the test value, then the failures were always of a brittle type. The mean test/prediction ratio for 24 tests is 1.12 with a standard deviation of 0.16. Considering the complexity of the problem, coupled with the fact that the tests involved a variety of failure modes, the consistency of the predictions is regarded as particularly good. A fundamental assumption of the design approach is the presumption of full plasticity within the connection. For specimens with low ratios of reinforcement, especially if only mesh of limited ductility was present and a

79

Composite connections in steel and concrete TABLE 1 Specimen Details of Flush Endplate Connections Used for the Validation

Source

Specimen

Najafi & Anderson Ref. 21

Testl Test3 Test6 Test7 Testl0 SCJ3 SCJ4 SCJ5 SCJ6 SCJ7 SCJ15 SJB10 SJB14 JX1 JX2 JY1 JC1 JC2 A2 A3 A4 C1 C2 C3 CJS-1 CJS-2 CJS-3 CJS-4 CJS-5 CJS-6

Xiao, Nethercot & Choo Ref. 5

Benussi, Smotlak & Zandonini Ref. 22 Law & Johnson Ref. :23

Aribert & Lachal Ref. 18

Li, Nethercot & Choo Ref. 16

Beam size

305 305 305 305 457 305 305 305 305 305 305

× 165UB40 x 165UB40 x 165UB40 x 165UB40 x 152UB52 x 165UB40 x 165UB40 × 165UB40 x 165UB40 x 165UB40 x 165UB40 IPE300 IPE300 457 x 191UB67 457 x 191UB67 457 × 191UB67 457 x 191UB67 457 × 191UB67 IPE360 HEA200 IPE360 IPE360 IPE360 IPE360 254 × 102UB25 254 × 102UB25 254 × 102UB25 254 × 102UB25 254 × 102UB25 254 × 102UB25

Column Slab section orientation bx d (mm) IN(F)* IN(F) IN(W) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F) IN(F)

Reinforcement ratio (%)

1100 x 120 1100 x 120 1100 x 120 1100 x 120 1100 x 120 1200 x 120 1200 x 120 1200 x 120 1200 x 120 1200 x 120 1200 × 120 1000 x 120 1000 x 120 1400 x 125 1400 x 125 1400 x 125 1400 x 125 1400 x 200 1000 x 120 1000 x 120 1000 x 120 1000 x 120 1000 × 120 1000 x 120 1000 x 110 1000 × 110 1000 × 110 1000 x 110 1000x 110 1000 x 110

1.0 0.5 0.2 1.5 1.0 0.2 1.0 1.0 1.2 1.0 1.0 0.71 1.21 0.72 0.72 0.72 0.72 0.45 0.53 0.53 0.53 0.74 0.74 0.74 0.90 0.90 0.90 0.90 0.90 0.90

* IN - - Internal joint. EX - - External joint. F - - Major axis. W - - Minor axis.

mesh fracture was the cause of the failure, this condition will not have been a t t a i n e d in e a c h o f t h e c o n n e c t i o n ' s c o m p o n e n t s . T h u s , t h e c r i t e r i a f o r t h e strength design as partial strength connections were not satisfied. Nor were criteria for design as semi-rigid connections satisfied because the rotation c a p a c i t i e s a r e t o o s m a l l a s d i s c u s s e d in R e f . 5. D e f o r m a t i o n c a p a c i t y o f c o m p o s i t e c o n n e c t i o n s w i l l b e d i s c u s s e d in a s e p a r a t e p a p e r . A t y p i c a l e x a m p l e o f t h i s is t e s t S C J 3 o f R e f . 5 f o r w h i c h t h e t e s t r e s u l t w a s o n l y 8 4 % o f t h i s

Y. Xiao, B. S. Choo, D. A. Nethercot

80

TABLE 2 Comparison Between Test and Calculated Connection Moment Capacities

Source

Najafi & Anderson Ref. 21

Xiao, Nethercot & Choo Re~ 5

Benussi, Smotlak &Zandoninin Re~ 22 Law & Johnson ReL 23

Aribea & Lachal* Re£ 18

Li, Nethercot& Choo Re~ 16

Total

Specimen

Beam properties

Connection properties

Ra~o (2)/(1)

Sagging moment (kNm)

Hogging moment (kNm)

Prediction (1) (kNm)

Test (2) (kNm)

Test 1 Test3 Test6 Test7 Test!0 SCJ3 SCJ4 SCJ5 SCJ6 SCJ7 SCJ 15 SJB 10 SJB 14

378 378 378 378 616 392 392 392 392 392 392 360 360

290 255 216 311 475 214 288 288 275 288 275 247 291

214 139 113 299 321 102 226 226 161 226 202 146 250

262 179 138 302 416 86 201 241 158 205 186 208 261

1.22 1.29 1.22 1.01 1.30 0.84 0.89 1.06 0.98 0.90 0.92 1.43 1.05

JX1 JX2 JY1 JCI JC2 A2 A3 A4 CI C2 C3 CJS-1 CJS-2 CJS-3 CJS-4 CJS-5 CJS-6

643 643 643 643 830 579 304 579 579 579 579 278 278 278 278 278 278

532 532 532 532 564 402 187 402 426 426 426 198 198 198 198 198 198

313 313 313 313 410 215 108 215 249 249 249 159 159 159 159 159 159

354 370 384 449 530 296 152 297 344 326 288 181 173 149 160 195 174

1.13 1.18 1.23 1.43 1.29 1.38 1.41 1.38 1.38 1.31 1.16 1.14 1.09 0.94 1.01 1.23 1.10

SD = 0.16

X = 14%

1.12

30

* Not included in the statistical summary.

81

Composite connections in steel and concrete

predicted value. However, such arrangements are not recommended - - more reinforcement should be used so as to produce a more ductile arrangement as well as to improve the moment capacity. For the arrangements that are most likely to be used in practice the predictions are generally slightly smaller than the equivale,nt test result. An assumption made when treating multiple bolt rows was that the connection' s neutral axis may be located at two-thirds of the vertical bolt pitch below the lowest row in tension. The sensitivity of this assumption has been examined using ~even tests taken from four different series. Figure 8 shows how the ratio of test to prediction varies as the separation point is moved from the conservative assumption of the top bolt row in compression (or = 1.0) to the last row in tension (or = 0). The average spread is from 1.36 to 1.05 and the proposed walue ((x = 1.67) gives a satisfactory 1.23. When only one row of bolts is in tension the neutral axis could be conservatively relocated to the second row of bolts. It is of interest that the sensitivity decreases with the (more usual) higher reinforcement ratios. Full studies of the separation ratio oz have been given in Ref. 15.

Partial depth endplates Fifteen tests from three research groups in UK, Italy and Canada [Van Dalen, K., Private Communication, 1993], 5"22 covering three different endplate 1.7 IP ----el----0.9P •

*

e~. 1.5 1.6

~

O.$P

*

0.7P 2/3P 0.6P

•

O.$P

~~..Upper b o u n d

1,4 ;~ 1.3 o,)

'

0.4P

1.2

--x--

0.3P

- - X ~ 0.2P O.IP

t

OP I% 0.9 SCJ5

I% ' Testl

0.5~% Test3

0.53% 0.74% n I A2 CI

0.90% 0.9

--

~

--

Ul~erbound

t - - - m-.

ClS-I

-

Used

Value

CJS-6 Lower

Fig. 8. Safety factors influenced by separation ratios.

bored

Y. Xiao, B. S. Choo, D. A. Nethercot

82

I

I

t Top

Bottom

Middle

(b)

(a)

(c)

Fig. 9. Different positions of partial depth endplate on the beam end.

locations (Fig. 9), have been used for validation purposes; these are listed in Table 3. Three of these tests 5 involved the use of endplates welded to the upper part of the beam. However, in the case of the cantilever specimen SCJ19 failure of the reinforcement anchorage behind the column limited the moment achieved. Thus, only two tests, in both of which failure was triggered by buckling of the beam web, have actually been used. TABLE 3 Specimen Details of Partial Depth Endplate Connections Used for the Validation

Source

Specimen

Beam size

Xiao, Nethercot & Choo (Top) Ref. 5 Van Dalen (Mid) Ref. 24

SCJ8 SCJ14 SCJI9 1 2 3 5 SCJ9 SJAIO SJA 14 SJAI4/I SJAI4/2 SJAI4/4 SCJIO SCJI3

305 x 165UB40 305 x 165UB40 305 x 165UB40 W360 x 39 W360 x 39 W360 x 39 W360 x 39 305 x 165UB40 IPE300 IPE300 IPE300 IPE300 IPE300 305 x 165UB40 305 x 165UB40

Xiao (Mid) Ref. 5 Zandonini & Benussi (Bot) Ref. 22

Xiao, Nethercot & Choo (BoO Ref. 5

Column Slab section Reinforcement orientation b ×d ratio (%) (ram) IN(F) IN(W) EX(W) IN(W) IN(W) IN(W) IN(W) IN(F) IN(W) IN(W) IN(W) IN(W) IN(F) IN(F) IN(W)

1200 x 120 1200 × 120 1200 x 120 2125 x 65 2125 x 65 2125 x 65 2125 x 65 1200 x 120 1000 x 120 I (X)0 x 120 1000 x 120 I(~0 x 120 I(X)O x 120 1200 × 120 1200 x 120

0.80 0.80 0.80 0.46 0.17 0.46 0.17 0.80 0.71 1.2 I 1.21 1.21 1.21 0.80 0.80

83

Composite connections in steel and concrete

In the case of endplates positioned at mid-depth of the beam four of the five available results are from tests on beam to beam specimens conducted by Van Dalen [Private Communication, 1993]. However, since the two beams were connected to either side of the web of a girder and were then subject to symmetrical loading, the test arrangement is actually equivalent to that of a minor axis beam to column specimen. In all five tests failure occurred by beam web buckling. Seven results are available for the potentially most effective arrangement, in which the endplate w a s attached to the bottom of the beam. For the specimens with unstiffened columns, buckling of the column web in compression was the main cause of failure, whilst for those tests in which column stiffening was employed excessive deflection of the composite slab eventually led to unloading. Omitting test SCJ19, the mean test/prediction ratio for the remaining 14 tests is 1.07 with a standard deviation of 0.07. In only one test SCJ10 was the prediction on the unsafe side. Thus, the design approach is considered as being satisfactorily validated for the case of partial depth endplate connections. TABLE 4

Composite Partial Depth Endplate Source

Specimen

Xiao, Nethercot & Choo (Top) Ref. 5 Van Dalen (Mid) Ref. 24

SCJ8 SCJ14 SCJ19* 1 2 3 5 Xiao (Mid) Ref. 5 SCJ9 Zandonini SJA 10 & Benussi (Bot) SJA14 Ref. 22 SJA14/1 SJAI4/2 SJA14/4 Xiao, Nethercot SCJI0 & Choo R,ef.5 SCJ13 Total

Beam properties

Connection properties

Ratio (2)/(1)

Sagging moment (kNm)

Hogging moment (kNm)

Prediction (1) (kNm)

Test (2) (kNm)

391 391 391 365 362 361 365 391 360 360 388 388 388 382 382

283 283 283 280 247 286 247 284 241 271 290 290 290 264 264

79 79 79 109 128 126 128 101 156 199 206 206 206 175 175

84 90 61 123 147 134 144 107 165 221 222 218 208 148 181

1.06 1.14 0.77 1.13 1.15 1.06 1.13 !.06 1.06 1.11 1.07 1.06 1.0 0.85 1.04

SD = 0.07

X = 6.8%

1.07

15

*SCJI9 is not included in the summary.

84

Y. Xiao, B. S. Choo, D. A. Nethercot

Designer choice Composite connections contain a comparatively large number of different components, each with the potential to exert a greater or lesser influence on moment capacity. Despite the 150 tests on composite connections already undertaken, 6 many combinations of variables remain untested and it is clearly not practical to expect to cover variations to the whole range of parameters by testing. The design model developed and validated above provides an opportunity to investigate the extent to which different factors will affect connection strength. Examples will be given herein for the key parameters such as reinforcement ratio, slab depth, etc., to be changed systematically to examine their influence on connection moment capacity and to compare this with the hogging moment capacity of the equivalent beam calculated by Refs 7 and 8. A limited study of this type has been undertaken and the results used to extract important trends in behaviour that can serve as a guide to efficient design. A particular feature of this is the levels of connection moment capacity (expressed as a fraction of beam hogging moment capacity) that may sensibly be achieved. These values may be linked directly to the approach to semicontinuous composite frame design reported elsewhere 9 and used to identify the most appropriate combinations for the efficient use of partial strength composite connections.

Flush endplates: effect of reinforcement ratio and slab depth Varying the slab depth and amount of reinforcement clearly provides a direct influence on moment capacity through changes in the lever arm and the tensile force, respectively. The basic connection arrangement was selected as that used for SCJ55 and illustrated in Fig. 10. Slab depths of 100 and 120 mm were considered and the reinforcement ratio p was varied from 0.4 to 2.2%. The specimen had a 10 mm thick endplate with 20 mm diameter 8.8 grade bolts. Steel grade was 43. Figure 11 shows how the moment capacity of the connection and the beam itself (in hogging) increase with the addition of more reinforcement. Providing sufficient reinforcement is used (1.7% for the 120 mm slab and 2.1% for the 100 mm slab), then connection capacities equal to the beam capacity may be achieved. Clearly adding to the reinforcement has more effect on the connection since beam capacities are much larger for the lower percentages. The explanation for this follows from the models used for the prediction in each case. For small reinforcement percentages the neutral axis will be located well into the steel beam. Whereas that part of the steel beam in tension actually contributes to hogging moment capacity, for the connection all tension must

Composite connections in steel and concrete

Shear studs Metal decking CF46 I[

'

\

I1

II

Welded web stiffener

II

85

Reinforcement ratio

II /

~

II

slab width = 1200 mm

Fig. 10. Composite flush endplate connection (SCJ5).

375 350 I

,j~,8~s~

....,a-~'

325

275 25O moment(!20r~deeO) •

0

Connection

moment(lOOmmdeep) . Hogging

175 ~ , / 1L50 0.4

moment(IOOmmdeep) I

0.6

0.8

1

I

1.2 1.4 1.6 Reinforcement ratio(%)

i

1

1.8

2

2.2

Fig. 11. Influence of reinforcement ratio on moment resistance.

be transfe,rred by the rebars acting with any bolts in tension. Once the neutral axis moves into the concrete slab the two moment capacities will become equal. Similm" calculations have been performed for the two reinforcement ratios of 1.0 and 1.4% with the slab depth being continuously varied. The results

Y. Xiao, B. S. Choo, D. A. Nethercot

86

plotted in Fig. 12 show a similar though less distinct trend with beam and connection capacities equalising at 185 and 140 mm. Thus, to achieve a connection moment capacity approaching the beam's hogging moment capacity requires a high percentage of reinforcement and/or a deep concrete slab. Use of the partial strength approach 9 would therefore appear to be the more attractive option. Partial depth endplate: effect of reinforcement ratio and slab depth Equivalent calculations have been conducted for partial depth endplates, starting with connections based on the arrangement of specimen SCJ10 (Fig. 13). This specimen also had a 10 mm thick endplate with 20 mm diameter 8.8 grade bolts. Steel grade was 43. In this case the endplate was located at the bottom of the beam. Figures 14 and 15 present the findings. Whilst the general pattern is similar to that observed for the full depth endplate with connection capacity increasing faster with increased reinforcement and slab depth, one important difference is apparent. Both curves become nearly parallel (at p - 1.6% in Fig. 14 and at 175 mm in Fig. 15). Connection capacities in excess of 94% of the beam's hogging moment capacity cannot therefore be achieved. Moving the endplate to the centre of the beam's depth further reduces the scope for achieving connection moment capacities approaching the beam hogi

I

I

i

ConneaJon moment(1.0%)

o

Hoggingmongnt

550 500

I

•

i

(,.o~) E 450

mor~nt(I.4~)

"o •~-P'

t"

Connection

350

--X--

Houing

h

.I

momcnt(l.4%)

~,~

L ~ f

-

150

{

10o 60

7o

80

90

too

11o

t20

130

14o

150

z6o

170

IS0

Depth of concrete slab (nan)

Fig. 12. Influence o f slab depth on moment resistance.

190 20o

Composite connections in steel and concrete

~.~..~.

\

Mo,.,~,~II

\

II

II

II /

87

~o~oo~o p.~

~-~I"/--~r)~~/-~ T/--~T/--~T/-I

'loll

,,

Fig. 13. Composite partial depth endplate connection (SCJ10).

400 3~) •~, 3OO

151) o

1oo 5O 0 0.3

0.5

0.7

0.9

1.1

1.3

1.5

1.7

1.9

2.1

2.3

Reinforcement ratio(%)

Fig. 14. Influence of reinforcemen~ ratio on partial depth endplate (bottom).

ging moment capacity. Studies ~5 show the two curves to be nearly parallel even at low reinforcement percentages or slab depths. Thus, the ratio Mco,IMhog becomes rather insensitive to changes in either p or d, and as endplate location moves towards the top of the beam so the relative level of connection strength that may be achieved is reduced.

Y. Xiao, B. S. Choo, D. A. Nethercot

88 500 450 "-'400 3so

v_~..f

~2so

jr""-" jr"

J

E 0

J

~2oo

"

Connection moment

J-

Hogging mon~nt

f I

~

i[r

150 I00

90

100

110

120

130

140

150

160

170

180

190

200

210

Slab depth(mm) Fig. 15. Influence of slab depth on partial depth endplate (bottom).

CONCLUSIONS From the establishment of the design model, validation against the available test data and initial parametric study, a reliable method has been provided for the prediction of the moment capacity of endplate types of composite connection. This method relates directly to recent relevant codes and may be presented as a complete design procedure. The document24 produced by the authors is an example of this. The following conclusions can be drawn for the present study: (1) A mathematical model has been developed for the prediction of the resistance moment of composite connections. The method is based on the rigid-plastic theory using the simplified concept of 'stress-blocks' which is in line with the calculation procedures for hogging moment capacity of composite beams in major codes. (2) All the assumptions for the model for this method have been experimentally calibrated against the authors' own test programme as well as other test results to give simple and straightforward formulations. (3) This model is suitable for different types of endplate connections with different dimensions, material properties and multi-rows of bolts. It covers the connection strength governed by different failure modes, including the controlling strength for connections without proper column web stiffening. (4) A total of 45 test specimens, including flush endplate and partial depth

Composite connections in steel and concrete

(5)

(6)

89

endplate connections has been examined using the moment capacity formulae proposed in the paper. The predicted and test ultimate moment resistances compared very well. This method is safe and accurate. The parametric study indicates that connection moment capacity (Moon) can reach the composite beam hogging moment (Mhog) when the reintbrcement ratio and/or slab depth is increased sufficiently for composite flush endplate connections. Connection moment can never reach the hogging moment capacity of the beam for other types of more flexible composite connections. However, the Mco,/Mhog ratios can be adjusted to a reasonably high figure by different joint arrangements, reinforcement ratios and slab depths.

ACKNOWLEDGEMENTS The work reported herein forms part of a study into the inclusion of joint behaviour in the design of composite frame. Various parts of this study have been supported by BRE, SERC and DTI. The authors acknowledge the helpful discussions conducted with many people during the course of this work, especially those with Dr D. B. Moore and Dr N. Jarrett of BRE, Dr R. M. Lawson of SCI, Dr C. Gibbons of Ove Arup and Partners and Dr D. Anderson of the University of Warwick.

REFERENCES 1. Barnard, P. R., Innovations in composite floor systems. Paper presented at the Canadian Structural Engineering Conference, Canadian Steel Industries Construction Council, p. 13, 1970. 2. Nethercot, D. A. & Zandonini, R., Recent studies of the performance and design of semi-rigid composite connections. American Society of Civil Engineers Structures Congress, Atlanta, USA, April 1994. 3. Commission of the European Communities, Eurocode no. 3: design of steel structures, part 1, general rules and rules for building. DD ENV 1993-1-1: BSI, London, 1992. 4. BSI, BS5950, Structural Use of Steelwork in Building Part 1: Code of Practice for Design in Simple and Continuous Construction: Hot Rolled Sections, London, 1!)90. 5. Xiao, Y., Choo, B. S. & Nethercot, D. A., Composite connection in steel and concrete. I. Experimental behaviour of composite beam-column connections. J. Construct. Steel Res. 31 (1994) 3-30. 6. Xiao, Y. & Nethercot, D. A., Database for composite connections. COSTC1 Working Group, C1-WD4-94-1, Lausanne, Switzerland, December 1993, pp. 1-4. 7. BSI, BS5950, Structural Use of Steelwork in Building Part 3.1: Design in Composite Construction, London, 1990.

90

Y. Xiao, B. S. Choo, D. A. Nethercot

8. Commission of the European Communities, Eurocode no. 4, design of composite steel and concrete structures, part 1.1: general rules and rules for buildings, DD ENV 1994-1-1, BSI, London, 1994. 9. Nethercot, D. A., Semi-rigid action and the design of non-sway composite frames. Engineering Structures (under review). 10. Li, T. Q., Choo, B. S. & Nethercot, D. A., Determination of rotation capacity requirements for steel and composite beams. J. Construct Steel Res. 32 (1995) 303-32. 11. Nethercot, D. A., Li, T. Q. & Choo, B. S., Required rotations and available degrees of moment redistribution. J. Construct. Steel Res. (in press). 12. Xiao, Y., Choo, B. S. & Nethercot, D. A., Design of semi-rigid composite beamcolumn connections. Proceedings of Conference of Building the Future: Innovations in Design, Materials and Construction, Brighton, UK, April 1993, pp. 391-406. 13. Johnson, R. P. & Hope-Gill, M. C., Semi-rigid joints in composite frames. IABSE, Ninth Congress, Prelim. Report, Amsterdam, May 1972, pp. 133 44. 14. BCSA/SCI, Joints in Simple Construction, Volume 1: Design Methods, Second edition, The Steel Construction Institute, 1993. 15. Xiao, Y., Behaviour of composite connections in steel and concrete. Ph.D thesis, University of Nottingham, February 1994. 16. Li, T. Q., The analysis and ductility requirement of semi-rigid composite frames. Ph.D. thesis, University of Nottingham, July 1994. 17. Ahmed, B. & Nethercot, D. A., Numerical modelling for semi-rigid composite connections. Progress report No. 1, Department of Civil Engineering, University of Nottingham, 1995. 18. Aribert, J. M. & Lachal, A., Experimental investigation of composite connections in global interpretation. Proceedings of The European Community COST C1 Conference on Semi-Rigid Joints, Strasbourg, France, October 1992, pp. 158-69. 19. Xiao, Y. & Nethercot, D. A., A new approach for moment capacity prediction for steel joints. Proceedings of 5th International Conference on Steel Structures, Jakarta, December 1994, pp. 327-31. 20. Xiao, Y. & Nethercot, D. A., Engineering bases of design proposal for composite connections in steel and concrete: no. 2 composite flush endplate connection (third version). Eureka Cimsteel E-130 Project - - Connection Detailing and Design, Department of Civil Engineering, University o f Nottingham, July 1994. 21. Anderson, D. & Najafi, A., Performance of composite connections. Major Axis End Plat Joints, 31 (1994) 31-57. 22. Benussi, F., Puhali, R. & Zandonini, R., Semi-rigid joints in steel-concrete composite frames. Construzioni Metalliche, 5 (1989) 1-28. 23. Law, C. L. C., Semi-rigid joints for composite frames, Ph.D thesis, University of Warwick, 1981. 24. Xiao, Y. & Nethercot, D. A., Design proposal for composite connections in steel and concrete (1) strength design of composite partial depth endplate connection. Preliminary Report for UK Working Group of Composite Connections, Department of Civil Engineering, University of Nottingham, September 1993.